- Email: [email protected]
Modelling pulverized coal combustion using air and O2+ recirculated flue gas as oxidant
Applied Thermal Engineering 83 (2015) 1e7
Contents lists available at ScienceDirect
Applied Thermal Engineering journal homepage: www.elsevier.com/locate/apthermeng
Research paper
Modelling pulverized coal combustion using air and O2 þ recirculated flue gas as oxidant ^ndio Rebola*, Joa ~o L.T. Azevedo Ama Mechanical Engineering Department, T ecnico Lisboa, Lisbon, Portugal
h i g h l i g h t s The flow structure is modified by the operating conditions. The wall temperature has a great influence on the wall incident radiation fluxes. Flames with similar temperatures have similar wall incident radiation fluxes. The influence of the CO2 levels on the wall incident radiation fluxes is marginal.
a r t i c l e i n f o
a b s t r a c t
Article history: Received 18 September 2014 Accepted 7 March 2015 Available online 17 March 2015
This work uses a numerical model to analyse the flame structure and the radiations heat transfers of pulverized coal firing with Air and in oxyfuel conditions. The numerical model considers an axisymmetric approximation for the flow and coal simulation based on a Lagrangian method. The combustion in the gas phase is handled by the eddy dissipation model combined with global kinetics. The coal conversion is based on the CPD volatilisation model and the char conversion by a first order kinetic model. Radiation properties are calculated based on a spectral line-based weighted-sum-of-grey-gases (SLW) model that was found appropriate for this situation. The radiation properties of particles are considered through the use of approximate functions for the absorption and scattering efficiencies. Calculations are presented for tests carried out in the RWEn power's 0.5MWth combustion test facility (CTF) with a Russian coal fired in a burner with 3% O2 in the flue gases, for air firing and simulated recirculated flue gases (RFG) with oxygen injection. The predicted incident radiation heat fluxes to the furnace walls are compared with experimental data. The calculations indicate that the flow structure is modified by the operating conditions leading to similar results when using the higher recirculation ratio and air firing. The computed wall incident heat fluxes present greater dependency of the wall temperatures than the CO2 level or the recirculation ratio. © 2015 Elsevier Ltd. All rights reserved.
Keywords: CFD modelling Oxycoal combustion Recirculated flue gases Radiation fluxes
1. Introduction Oxycoal combustion is one of the technologies that is considered to control green gases emissions of the great coal power plants. This technology produces a high concentration of CO2 on the flue gases stream and has the potential to simplify the process of carbon capture and storage (CCS), avoiding the nitrogen separation step necessary on conventional combustion with air. In the
* Corresponding author. Mechanical Engineering Department, Instituto Superior cnico, Av. Rovisco Pais, 1049-001 Lisboa, Portugal. Te E-mail address: [email protected] (A. Rebola). http://dx.doi.org/10.1016/j.applthermaleng.2015.03.008 1359-4311/© 2015 Elsevier Ltd. All rights reserved.
literature several authors [1e3] presents reviews of this combustion process. Replacing the N2 by CO2 affects the coal particles combustion, single particles studies show that at low oxygen levels the presence of CO2 increases the ignition delay [4,5]. To obtain a similar ignition time as air, in a mixture of O2/CO2 it is necessary a superior level of oxygen [4]. After ignition, under conditions where combustion is controlled by chemical kinetics, the particle conversion in mixtures of O2/CO2 and O2/N2 appears to be similar [5]. However at higher temperatures, when external mass transfer influences the combustion process, faster conversion was found in N2 suggesting that the lower molecular diffusion coefficient of O2 in CO2 lower the char conversion rate [5]. On the other hand, the experimental result presented by Rathnam et al. [6] shows the potential of the CO2
2
A. Rebola, J.L.T. Azevedo / Applied Thermal Engineering 83 (2015) 1e7
gasification to increase the particle mass release during volatilization and to modify the morphology of the char particle. This work also demonstrated the importance of CO2 gasification reactions on char consumption in mediums with lows levels of O2 and high temperature. Other consequence of changing N2 by, CO2 and water vapour, in the oxidant stream, is the increase of the flue gases emissivity and changes on the heat transfer in the furnaces. Nevertheless, experimental results obtained in pilot scales furnaces show that radiation heat fluxes are mainly determined by the temperature [7]. In oxycoal combustion the in-furnace heat transfer and temperature profiles can be made similar to air firing condition by adjusting O2 concentration and recirculation ratio [8]. However, to obtain a similar flame temperature as in air firing situation, the different heat capacity and densities of the CO2 relatively to N2, will change the mass flows and velocities in the primary and secondary flows and as a consequence the burners aerodynamics, which result in changes in the flame shape, fuel ignition properties, flame propagation, and residence time of the coal particles [9]. Changes to the burner dimension and/or increase the swirl may be required to maintain a similar flame as air in oxycoal firing conditions [9]. Oxycoal combustion was also study using numerical simulations tools, at the Stuttgard University [10] a global homogeneous combustion model that accounts for the chemical effects of a high CO2 concentration under oxyfuel conditions by means of reversible reaction pathways was developed and coupled with turbulence by the EDC model. The model was used to simulate the combustion of dried lignite using air and a mixture of O2/CO2 as oxidant. The char conversion model was derived from a general physical basis to account for boundary layer diffusion and pore diffusion with chemical reaction. Kinetic constants for char gasification reactions with oxygen, carbon dioxide and water vapour were considered. Despite discrepancies in the vicinity of the burner, there is overall satisfying agreement between measured and predicted values. Referring to the O2 concentration profile, the model seems to predict the ignition of the flame too late. Dobrin et al. [11] used a model where the homogeneous chemical reaction was modelled by the Eddy Dissipation Model/Finite Rate EDM/FR approach. The devolatilisation process was modelled using the CPD model and the char burnout rate was obtained from apparent surface kinetic and diffusion mass transfer rates. The model provides a good representation of the velocity, oxygen and temperature distributions. Nevertheless, the predictions present larger gradients than experimental data. A similar model was used to improve the burner design [12] and obtain stable flames under oxyfuel conditions reducing the oxygen volume fraction on the burner stream and to investigate a multi-burner utility scale furnace operating in both air and oxyfuel combustion conditions [13]. This paper presents the application of a numerical model to the RWEn power's 0.5 MWth combustion test facility (CTF) for coal flames in air and mixtures of CO2 and O2 to simulated recirculated flue gases (RFG) [14]. Particular attention was given to the radiation heat transfer. The computed radiation heat fluxes at wall are compared with experimental data for air and oxycoal firing conditions with recirculations ratio between 0.65 and 0.75.
the turbulence/chemistry interaction. In the EDM it is assumed that the reaction rate of a species is fully controlled by turbulent mixing. The EDM/FR calculates the reaction rate of a species as the minimum between the mixing rate and a kinetics rate, which is evaluated from an Arrhenius equation based on the mean properties. The species considered as the result of the devolatilisation process are considered to react in a two step irreversible reactions, in the first step the fuel is converted to form CO and water and the second step is the CO oxidation to CO2. H2 is assumed to convert immediately to H2O. CO is also present in the volatiles and is formed from the char combustion, so it enters directly to the second step. Table 1 presents the parameters of the global reaction rates used in the simulations obtained from Refs. [16,17] and defined as:
Ea ½Fuela ½O2 b ½H2 Oc Rch ¼ A exp RT
h . i kmol m3 s
(1)
The coal is simulated in a Lagrangian framework tracking representative particles within the furnace with reflection at the walls until they exit the furnace. The energy balance to the particle is used to calculate its temperature and the particles in sequence go through drying, devolatilisation and char conversion. Drying is assumed to occur at 100 C requiring heat from the surroundings, while during char conversion the energy released heats the particle. The heat transfer between the particle and the continuum medium occurs by convection and radiation. The devolatilisation process is modelled using the Chemical Percolation Devolatilisation (CPD) model, where the detailed mechanisms of devolatilisation reactions, including bridge breaking and rearranging, light gas release, tar evaporation, and cross-linking, are considered quantitatively. Therefore, the model predicts not only the rate of volatile release, the yields of the light gas, tar (CxHyOz) and char but also the light gas species distribution in volatiles, namely (CH4, CO, CO2, H2O and other light gases represented by C2H4). In the present work, the model prediction of volatile species was coupled with chemical kinetics for the gas species combustion as mentioned before. For the char conversion the oxidation reaction producing CO due to the small particle sizes and large temperature the oxidation reaction was considered first order with the pre-exponential constant A ¼ 0.005 kg/sm2Pa and activation energy 74,000 kJ/kmol. The thermal radiation is calculated using the discrete ordinates method. The radiative properties of the participating medium were modelled by the weighted-sum-of-grey gases model (SLW) [18] and the radiation properties of particles are calculated from approximations to the Mie theory [19]. The model considers distinct approximate functions to calculate the absorption and scattering coefficients of coal, char and ash particles.
3. Test cases The conditions considered for the simulations were those from the tests made at RWEn power's 0.5 MWth combustion test facility (CTF) [14]. As shown in Fig. 1 the CTF has a refractory lined combustion chamber composed by a block with 0.8 0.8 m inner cross-
2. Numerical model The model is based on the numerical solution of the Favreaveraged balance equations for mass of individual species, momentum, turbulence quantities and energy. The SIMPLER algorithm was used for velocity-pressure coupling and the standard k-ε model was used to model the turbulence. Combustion in the gas phase is calculated based on the Eddy Dissipation Model [15]/Finite Rate EDM/FR approach to describe
Table 1 Parameters for global reaction rates (*modelled as a long chain hydrocarbon with molecular weight of 120 kg/kmol). Fuel
A
CH4 C2H4 Tar* CO
5.01 2.42 1.98 2.24
1011 1010 1010 1012
Ea/R
a
b
c
Ref.
24,334 15,083 15,083 20,110
0.7 0.10 0.25 1.00
0.8 1.65 1.50 0.25
0 0 0 0.5
[16] [17] [17] [16]
A. Rebola, J.L.T. Azevedo / Applied Thermal Engineering 83 (2015) 1e7
3
Fig. 1. Schematic representation of the CTF [14].
section and 4 m long and a final convergent block with 2.18 m length and an end cross-section of 0.3 0.3 m. The installation is equipped with a 0.5 MWth (IFRF) aerodynamically air staged burner, a schematic representations of the burner is presented in Fig. 2. The CTF was originally designed for air-firing and was retrofitted to once-through oxyfuel system and the conditions considered in the present paper are those for a Russian coal fired with 3% O2 in the flue gases simulating dry recycling. The recirculation ratio, RR, is defined as:
RR ¼
mRFG mRFG þ mPFG
(2)
mRFG is the recirculated gas mass flow rate and mPFG is the flue gas mass flow rate. Table 2 shows the test conditions for the case of air firing and oxycoal combustion for the four recirculation ratios that were considered for the simulations. The temperatures of the primary and secondary stream were 70 and 270 C, respectively. The O2 concentration in the primary stream was maintained at 21% (V/V) for all recycle ratios and the primary stream velocity was maintained constant. Table 3 presents the composition and calorific
value of the Russian coal. The particle size distribution was approximated by a Rossin Rammeler distribution based on the diameters distributions of the Russian coal presented in Ref. [20] and shown in Table 4. The calculations were carried out using a staggered, two dimensional axisymmetric, non-uniform grid. The grid had a higher density near the burner region and in the vicinity of the centreline containing 126 120 cells and the computational domain extends over 7822 451 mm (axial radial direction). The radial dimension was obtained maintaining the cross section area of the axisymmetric mesh equal to the real furnace, and the axial dimension of the grid guarantees that the exit boundary is more than ten diameter away of the last cross area change. The boundary at the burner axis was treated as axisymmetric, and the downstream boundary was treated as outflow. At the burner exit the radial velocity component was taken as zero, the secondary stream tangential velocity was estimate based on the swirl number and the axial velocity was calculated from mass flow rate of the transport and secondary fluid presented in Table 2. The inlet turbulent kinetic energy, kin, was estimated assuming that the inlet turbulence intensity is 10%. The inlet values of the dissipation
Fig. 2. Schematic representation of the burner [14].
4
A. Rebola, J.L.T. Azevedo / Applied Thermal Engineering 83 (2015) 1e7
Table 2 Test conditions (*N2 mass flow). Recirculation ratio [%]
Coal [kg/h]
Primary fluid O2 [kg/h]
Air RR 65 RR 68 RR 72 RR 75
68.8 68.8 68.8 68.8 68.8
25 25 25 25 25
Secondary fluid CO2 [kg/h] 83 130 130 130 130
*
[%] mass
Moisture Volatile matter Ash Fixed carbon Carbon Hydrogen Nitrogen Sulphur Oxygen Gross calorific value [kJ/kg]
6.23 33.55 11.98 48.27 65.91 4.59 2.09 0.34 8.86 27,098
h h i i 2pklðTw T∞ Þ 4 G εsTw ð1 εÞG þ hðTin Tw Þ Aint ¼ lnðRext =Rint Þ (3) G represents the incident heat flux at the wall, Tw the interior wall temperature, T∞ the exterior wall temperature considered 298 K, h the convection heat transfer coefficient, Ain the interior wall area, l the control volume length near the wall, s the StefaneBoltzmann constant, k the wall thermal conductivity, ε the wall emissivity, Tin the temperature of the inside medium at the wall,
Diameters [mm]
[%] mass
300 212 150 106 75
99.7 99.2 97.6 91.6 77.8
141 129 130 132 134
CO2 [kg/h] 463 226 278 364 448
*
O2 [%] v/v 21 44 39 33 29
4. Numerical results and discussion
rate of turbulent kinetic energy were determined from: εin ¼ C kin/l. The mixing length, l, was taken as 5% of the inlet area square root the constant C was set equal to 0.09. Unfortunately the CTF wall temperature and emissivity are not available. To perform the simulations the internal wall temperature distribution of each case was estimated based on the energy balance at the wall, assuming that the heat flux thought the refractory wall is equal to the radiation flux absorbed by the wall at each section of the furnace. The furnace outside wall temperature was considered 298 K, the values 0.3 m and 2.74 W/(mK) was considered to the wall thickness and the thermal conductivity of the refractory respectively and for the wall emissivity the value of 0.85 was considered. The wall temperature distributions obtained by this approach for each case are presented below. We also perform simulations for the Air firing conditions with different wall properties to evaluate the influence of the wall boundary conditions on the results. On these simulations the wall temperature was calculated inside the code based on the local energy balance at the wall, expressed as:
Table 4 Coal particle size distribution (experimental values) [20].
21 21 21 21 21
O2 [kg/h]
and Rint and Rext the interior and exterior radius of the wall respectively. Simulations A and B were performed with a wall thermal conductivity of 3.54 W/mK and a wall emissivity of 0.85 and 0.5 respectively. Simulation C was performed with a wall emissivity of 0.85 and a lager thermal conductivity 4.2 W/mK.
Table 3 Proximate and ultimate analyses of coal as received and averaged [14]. Russian coal
O2 [%] v/v
Fig. 3 presents the stream lines and coal particle consumption rate, predicted by the model, for the five study cases s. The predicted temperature and oxygen volumetric fractions distributions are presented in Fig. 4. For all cases the flow, in the near burner region, is characterized by a short internal recirculation zone, attached to the quarl, and a large external recirculation zone. The model predicts ignitions, for the five cases, inside the internal recirculations zone, near the quarl. As it is showed by the low level of oxygen and the relatively high temperatures predicted in this zone. The higher temperature computed inside the internal recirculation zone for the Air firing conditions comparatively to the cases with recirculation could be explained by the low mass flow rate, for the same amount of oxygen, of the burner primary stream and the lower heat capacity of the N2 comparatively to CO2. As shown in Table 2 for the recirculation cases to keep the same oxygen flow rate and volumetric fraction in the primary stream, as the Air case, it is necessary to increase the burner primary mass flow rate. Further due to the reduction of the secondary flow rate for the lowers recirculation ratio, the primary to secondary momentum ratio increases and therefore the effect of the swirled secondary stream is smaller leading to a less intense internal and external recirculation zones. These changes, in the near burner zone flow, have a significant influence on the flame structure. For the Air firing conditions and the higher recirculation rates the particle consumption occurs essentially inside the external recirculation zone and in the shear layer between the main jet and the external recirculation zone. For the lowers recirculation ratio, the flow capacity to drag the particles into the external recirculation zone is reduced, as a consequence, the conversion of the coal particles occurs essentially in the shear layer between the main jet and the external recirculation zone. The structure of the flame approximates to a jet type flame for the lowers recirculation ratio. This is corroborated by the computed distributions of oxygen and temperature. For the lower recirculation ratio the oxygen level are very high in the external recirculation zone, as a consequence of the high levels of oxygen in the secondary stream and the lower penetration of the particles in the external recirculation zone. These cases present high temperatures and low levels of oxygen in the shear layer between the main jet and the external recirculation zone, where essentially occurs the combustion process. The RR 75 is the case with the oxygen level and mass flow rate of the secondary stream closer to the Air firing conditions. The flow, oxygen distributions and coal particles mass release computed for these two cases have more similarities comparing to the others cases. These results clearly indicate that, for combustion with
A. Rebola, J.L.T. Azevedo / Applied Thermal Engineering 83 (2015) 1e7
5
Fig. 3. Stream lines and particle consumption distribution. (Wall temperatures presented in Fig. 5, wall emissivity 0.85).
recirculated flue gases to obtain flames with the same structure as the Air firing conditions it is necessary to maintain the burner aerodynamic and have the appropriate level of oxygen in the burner stream. The computed temperatures increase with the reductions of the recirculation ratio due to the reduction of the mass flow rate of gases, for the same thermal input from coal, as the recirculation ration decreases. Despite the higher recirculation ratios present an oxygen distribution and flow structure more similar to the Air firing conditions, the RR 72 is the case with the computed temperature distribution closer to the Air case. The lower temperatures predicted for the RR 75, comparatively to the Air case, are explained by the higher heat capacity of CO2 comparatively to N2, these two cases have similar burner mass fluxes. For the RR 72 the difference of the heat capacity between CO2 and N2 is compensated with the reduction of the CO2 mass flow rate comparatively to the N2 of the Air firing conditions.
Fig. 5 presents the wall temperature distribution used in the simulations obtained based on the experimental incident wall flux and the comparison between the experimental results and the computed heat wall incident fluxes, for the five cases studies. It can be observed that in general there is good agreement between the computational results and the experimental data. Nevertheless the computed wall incident heat fluxes over predict the measured values for all cases. The RR 72 is the case with the computed wall incident closer to the Air firing conditions. These two cases have also a similar computed temperature distributions as shown in Fig. 4. Fig. 6 presents a comparison between the calculated incident heat fluxes obtained for Air, RR 65 and RR 75, considering the prescribed wall temperature of the Air firing condition, presented in Fig. 5. The difference between the incident heat fluxes, for the different cases, is much lower than the previous results shown in Fig. 5. These results show the influence of the wall temperature on
Fig. 4. Temperature and O2 wet volume distribution. (Wall temperatures presented in Fig. 5, wall emissivity 0.85).
6
A. Rebola, J.L.T. Azevedo / Applied Thermal Engineering 83 (2015) 1e7
mass flow rate of the RR 75 comparatively to the Air case. An explanation for the results presented in Fig. 6 is the higher contribution of the medium to radiation flux, due the higher levels of CO2, on the RR 75 comparatively to the Air case. Fig. 7 presents the computed wall temperature, incident and absorbed radiation heat fluxes of simulations A, B and C, as shows the comparison between the results of simulation A and B, a significant change of the wall emissivity had a marginal effect on the results. The incident heat flux is part from the emitted and reflected walls and from the participation medium gas and particles. A decrease of the wall emissivity causes a reduction of the heat flux emitted by the walls and an increase of the reflected flux, the results suggest that the wall effects cancels each others, as a consequence changes of the wall emissivity value have a small effect on the computed results. The reduction of the wall emissivity causes
Fig. 5. Prescribed wall temperature and predicted wall incident heat flux. (Wall emissivity 0.85).
the computed wall incident heat fluxes. The higher fluxes are obtained for the RR 65 firing conditions, as explained before, the lower flow rate of gases for the same heat input from the coal will provide higher temperature and as a consequence higher fluxes. For the RR 75 conditions the predicted wall fluxes are slightly higher than the case of Air firing on the early section of the CTF. The predicted temperatures, in the near burner zone, of the RR 75 are slight lowers than the Air case, a consequence of the higher calorific value of the CO2 comparatively to the N2 and the superior burner
Fig. 6. Wall incident radiation flux for Air, RR 65 and RR 75 with the prescribed temperature of the Air case presented in Fig. 5. (Wall emissivity 0.85).
Fig. 7. Predicted wall temperature, incident and absorbed radiation heat flux for different wall properties. (Simulations A, B and C).
A. Rebola, J.L.T. Azevedo / Applied Thermal Engineering 83 (2015) 1e7
just a small reduction on the computed wall temperature and the radiated heat flux absorbed by the walls in the early section of the CTF. On the other hand, changes of the wall thermal conductivity have a larger influence on the results, particularly on the computed wall incident heat flux. The comparison between the results of simulation A and C shows, as expected, an increase of the computed wall temperature with the reduction of the considered wall thermal conductivity. As already explained, the higher temperatures of the wall and consequently of the interior medium increase the wall incident flux. The results also show that, a small increase of the computed wall temperature, of approximately 3%, increases the computed incident heat flux in more than 10% and reduces the absorbed wall flux in about 7%. These results clearly show the importance of the wall temperature especially on the accuracy of the computed incident heat flux at the wall. The direct comparison of incident heat fluxes without analysing the effect of wall temperature may provide misleading conclusions. This is not the case for water cooled membrane walls in boilers where the thermal resistance of the wall is much smaller and the wall temperature much lower. However for hot refractory wall test furnaces where there is a high dependency between the internal wall temperature and the incident heat flux, care should be given to the interpretation of the results. 5. Conclusions This work presents a simulation of coal combustion using air and mixtures of CO2 and oxygen as comburent, giving special attentions to the radiations wall incident heat fluxes and the flames structure. The computational results are compared with experimental data obtained in a 500 kWth combustion test facility at RWEn. The numerical model considers an axisymmetric approximation for the flow and coal simulation based on a Lagrangian method. The calculations indicate that the flame structure is modified by the operating conditions, with the reduction of the recirculation ratio due to the reduction of the secondary flow rate, the effect of the swirled secondary stream becomes smaller and the structure of the flame approximates to a jet type flame. The computed results also show that oxy-flames with a level of oxygen of 29% volume on the secondary burner stream and with a similar momentum as the Air firing conditions present a flame structure similar to the Air case flame. With the appropriate wall temperature, the radiation heat fluxes predicted by the model are in agreement and have the same trends of the experimental data. The numerical result shows that flames with near computed temperature distribution presents similar computed wall incident heat radiation fluxes. The numerical results also show that the computed wall incident heat fluxes are much more influenced by the wall temperatures than by the recirculation ratio or the CO2 levels. Therefore the analysis of the radiation fluxes obtained in hot refractory wall test furnaces should be done with care to avoid misleading conclusions.
7
Acknowledgements ^ncia e a TecThis work was supported by Fundaç~ ao para a Cie nologia (FCT), through IDMEC, under LAETA Pest-OE/EME/LA0022 and the project BOFCom financially supported by the Research Fund for Coal and Steel under contract RFCReCTe2006e00010.
References [1] T. Wall, Y. Liu, C. Spero, L. Elliott, S. Khare, R. Rathnam, F. Zeenathal, B. Moghtaderi, B. Buhre, C. Sheng, R. Gupta, T. Yamada, K. Makino, J. Yu, An overview on oxyfuel coal combustion-state of the art research and technology development, Chem. Eng. Res. Des. 87 (2009) 1003e1016. [2] G. Scheffknecht, L. Al-Makhadmed, U. Schnell, J. Maier, Oxy-fuel coal combustion-A review of the current state-of-the-art, Int. J. Greenh. Gas Control 5 (2011) S16eS35. [3] L. Cheng, S.Z. Yong, A.F. Ghoniem, Oxy-fuel combustion of pulverized coal: characterization, fundamentals, stabilization and CFD modelling, Prog. Energy Combust. Sci. 38 (2012) 156e214. [4] A. Molina, C.R. Shaddix, Ignition and devolatilisation of pulverized bituminous coal particles during oxygen/carbon dioxide coal combustion, Proc. Combust. Inst. 31 (2007) 1905e1912. [5] J. Brix, P.A. Jensen, A.D. Jensen, Coal devolatilisation and char conversion under suspension fired conditions in O2/N2 and O2/CO2 atmospheres, Fuel 89 (2010) 3373e3380. [6] R.K. Rathnam, L.K. Elliot, T.F. Wall, Y. Liu, B. Moghtaderi, Differences in reactivity of pulverised coal in air (O2/N2) and oxy-fuel (O2/CO2) conditions, Fuel Process. Technol. 90 (2009) 797e802. [7] K. Anderson, R. Johansson, S. Hj€ artstam, F. Johnsson, B. Leckner, Radiation intensity of lignite-fired oxy-fuel flames, Exp. Therm. Fluid Sci. 33 (2008) 67e76. [8] Y. Tan, E. Croiset, M.A. Douglas, K.V. Thambimuthu, Combustion characteristics of coal in a mixture of oxygen and recycled flue gas, Fuel 85 (2006) 507e512. [9] S.P. Khare, T.F. Wall, A.Z. Faruda, Y. Liu, B. Moghtaderi, R.P. Gupta, Factors influencing the ignition of flames from air-fired swirl pf burners retrofitted to oxy-fuel, Fuel 67 (2008) 1042e1049. [10] S. Leiser, Numerical Simulation of Oxy-fuel Combustion (Ph.D. thesis), University of Stuttgart, Germany, Shaker-Verlag Aachen, 2011, ISBN 978-3-83229722-0. [11] D. Toporov, P. Bocian, P. Heil, A. Kellermann, H. Stadler, S. Tschunko, € rster, R. Kneer, Detailed investigation of a pulverized fuel swirl flame in M. Fo CO2/O2 atmospheres, Combust. Flame 155 (2008) 605e618. € rster, R. Kneer, Development [12] P. Heil, D. Toporov, H. Stadler, S. Tschunko, M. Fo of an oxycoal swirl burner operating at low O2 concentrations, Fuel 88 (2009) 1269e1274. €ster, R. Kneer, Experimental and [13] M. Habermehl, J. Erfurth, D. Toporov, M. Fo numerical investigation on a swirl oxycoal flame, Appl. Therm. Eng. 49 (2012) 161e169. [14] J.P. Smart, P. O'Nions, G.S. Riley, Radiation and convective heat transfer and, burnout in oxy-coal combustion, Fuel 89 (2010) 2468e2476. [15] B.F. Magnussen, B.H. Hjertager, On mathematical modeling of turbulent combustion with special emphasis on soot formation and combustion, Proc. Combust. Inst. 16 (1977) 719e729. [16] F.L. Dryer, I. Glassman, High-temperature oxidation of CO and CH4, Proc. Combust. Inst. 14 (1973) 987e1003. [17] C.K. Westbrook, F.L. Dryer, Simplified reaction mechanisms for the oxidation of hydrocarbon fuels in flames, Combust. Sci. Technol. 27 (1981) 31e43. [18] M.K. Denison, A Spectral Line-based Weighted-sum-of-gray-gases Model for Arbitrary RTE Solvers (PhD thesis), Bringham Young University, United States of America, 1994. [19] W.A. Fiveland, W.J. Oberjohn, D.K. Cornelius, A Numerical Model for Predicting Furnace Performance in Axisymmetric Geometries, Department of Energy, United States of America, 1985. Report DOE/PC/40265e9. [20] J.P. Smart, R. Patel, G.S. Riley, Oxy-fuel combustion of coal and biomass, the effect on radiative and convective heat transfer and burnout, Combust. Flame 175 (2010) 2230e2240.
Contents lists available at ScienceDirect
Applied Thermal Engineering journal homepage: www.elsevier.com/locate/apthermeng
Research paper
Modelling pulverized coal combustion using air and O2 þ recirculated flue gas as oxidant ^ndio Rebola*, Joa ~o L.T. Azevedo Ama Mechanical Engineering Department, T ecnico Lisboa, Lisbon, Portugal
h i g h l i g h t s The flow structure is modified by the operating conditions. The wall temperature has a great influence on the wall incident radiation fluxes. Flames with similar temperatures have similar wall incident radiation fluxes. The influence of the CO2 levels on the wall incident radiation fluxes is marginal.
a r t i c l e i n f o
a b s t r a c t
Article history: Received 18 September 2014 Accepted 7 March 2015 Available online 17 March 2015
This work uses a numerical model to analyse the flame structure and the radiations heat transfers of pulverized coal firing with Air and in oxyfuel conditions. The numerical model considers an axisymmetric approximation for the flow and coal simulation based on a Lagrangian method. The combustion in the gas phase is handled by the eddy dissipation model combined with global kinetics. The coal conversion is based on the CPD volatilisation model and the char conversion by a first order kinetic model. Radiation properties are calculated based on a spectral line-based weighted-sum-of-grey-gases (SLW) model that was found appropriate for this situation. The radiation properties of particles are considered through the use of approximate functions for the absorption and scattering efficiencies. Calculations are presented for tests carried out in the RWEn power's 0.5MWth combustion test facility (CTF) with a Russian coal fired in a burner with 3% O2 in the flue gases, for air firing and simulated recirculated flue gases (RFG) with oxygen injection. The predicted incident radiation heat fluxes to the furnace walls are compared with experimental data. The calculations indicate that the flow structure is modified by the operating conditions leading to similar results when using the higher recirculation ratio and air firing. The computed wall incident heat fluxes present greater dependency of the wall temperatures than the CO2 level or the recirculation ratio. © 2015 Elsevier Ltd. All rights reserved.
Keywords: CFD modelling Oxycoal combustion Recirculated flue gases Radiation fluxes
1. Introduction Oxycoal combustion is one of the technologies that is considered to control green gases emissions of the great coal power plants. This technology produces a high concentration of CO2 on the flue gases stream and has the potential to simplify the process of carbon capture and storage (CCS), avoiding the nitrogen separation step necessary on conventional combustion with air. In the
* Corresponding author. Mechanical Engineering Department, Instituto Superior cnico, Av. Rovisco Pais, 1049-001 Lisboa, Portugal. Te E-mail address: [email protected] (A. Rebola). http://dx.doi.org/10.1016/j.applthermaleng.2015.03.008 1359-4311/© 2015 Elsevier Ltd. All rights reserved.
literature several authors [1e3] presents reviews of this combustion process. Replacing the N2 by CO2 affects the coal particles combustion, single particles studies show that at low oxygen levels the presence of CO2 increases the ignition delay [4,5]. To obtain a similar ignition time as air, in a mixture of O2/CO2 it is necessary a superior level of oxygen [4]. After ignition, under conditions where combustion is controlled by chemical kinetics, the particle conversion in mixtures of O2/CO2 and O2/N2 appears to be similar [5]. However at higher temperatures, when external mass transfer influences the combustion process, faster conversion was found in N2 suggesting that the lower molecular diffusion coefficient of O2 in CO2 lower the char conversion rate [5]. On the other hand, the experimental result presented by Rathnam et al. [6] shows the potential of the CO2
2
A. Rebola, J.L.T. Azevedo / Applied Thermal Engineering 83 (2015) 1e7
gasification to increase the particle mass release during volatilization and to modify the morphology of the char particle. This work also demonstrated the importance of CO2 gasification reactions on char consumption in mediums with lows levels of O2 and high temperature. Other consequence of changing N2 by, CO2 and water vapour, in the oxidant stream, is the increase of the flue gases emissivity and changes on the heat transfer in the furnaces. Nevertheless, experimental results obtained in pilot scales furnaces show that radiation heat fluxes are mainly determined by the temperature [7]. In oxycoal combustion the in-furnace heat transfer and temperature profiles can be made similar to air firing condition by adjusting O2 concentration and recirculation ratio [8]. However, to obtain a similar flame temperature as in air firing situation, the different heat capacity and densities of the CO2 relatively to N2, will change the mass flows and velocities in the primary and secondary flows and as a consequence the burners aerodynamics, which result in changes in the flame shape, fuel ignition properties, flame propagation, and residence time of the coal particles [9]. Changes to the burner dimension and/or increase the swirl may be required to maintain a similar flame as air in oxycoal firing conditions [9]. Oxycoal combustion was also study using numerical simulations tools, at the Stuttgard University [10] a global homogeneous combustion model that accounts for the chemical effects of a high CO2 concentration under oxyfuel conditions by means of reversible reaction pathways was developed and coupled with turbulence by the EDC model. The model was used to simulate the combustion of dried lignite using air and a mixture of O2/CO2 as oxidant. The char conversion model was derived from a general physical basis to account for boundary layer diffusion and pore diffusion with chemical reaction. Kinetic constants for char gasification reactions with oxygen, carbon dioxide and water vapour were considered. Despite discrepancies in the vicinity of the burner, there is overall satisfying agreement between measured and predicted values. Referring to the O2 concentration profile, the model seems to predict the ignition of the flame too late. Dobrin et al. [11] used a model where the homogeneous chemical reaction was modelled by the Eddy Dissipation Model/Finite Rate EDM/FR approach. The devolatilisation process was modelled using the CPD model and the char burnout rate was obtained from apparent surface kinetic and diffusion mass transfer rates. The model provides a good representation of the velocity, oxygen and temperature distributions. Nevertheless, the predictions present larger gradients than experimental data. A similar model was used to improve the burner design [12] and obtain stable flames under oxyfuel conditions reducing the oxygen volume fraction on the burner stream and to investigate a multi-burner utility scale furnace operating in both air and oxyfuel combustion conditions [13]. This paper presents the application of a numerical model to the RWEn power's 0.5 MWth combustion test facility (CTF) for coal flames in air and mixtures of CO2 and O2 to simulated recirculated flue gases (RFG) [14]. Particular attention was given to the radiation heat transfer. The computed radiation heat fluxes at wall are compared with experimental data for air and oxycoal firing conditions with recirculations ratio between 0.65 and 0.75.
the turbulence/chemistry interaction. In the EDM it is assumed that the reaction rate of a species is fully controlled by turbulent mixing. The EDM/FR calculates the reaction rate of a species as the minimum between the mixing rate and a kinetics rate, which is evaluated from an Arrhenius equation based on the mean properties. The species considered as the result of the devolatilisation process are considered to react in a two step irreversible reactions, in the first step the fuel is converted to form CO and water and the second step is the CO oxidation to CO2. H2 is assumed to convert immediately to H2O. CO is also present in the volatiles and is formed from the char combustion, so it enters directly to the second step. Table 1 presents the parameters of the global reaction rates used in the simulations obtained from Refs. [16,17] and defined as:
Ea ½Fuela ½O2 b ½H2 Oc Rch ¼ A exp RT
h . i kmol m3 s
(1)
The coal is simulated in a Lagrangian framework tracking representative particles within the furnace with reflection at the walls until they exit the furnace. The energy balance to the particle is used to calculate its temperature and the particles in sequence go through drying, devolatilisation and char conversion. Drying is assumed to occur at 100 C requiring heat from the surroundings, while during char conversion the energy released heats the particle. The heat transfer between the particle and the continuum medium occurs by convection and radiation. The devolatilisation process is modelled using the Chemical Percolation Devolatilisation (CPD) model, where the detailed mechanisms of devolatilisation reactions, including bridge breaking and rearranging, light gas release, tar evaporation, and cross-linking, are considered quantitatively. Therefore, the model predicts not only the rate of volatile release, the yields of the light gas, tar (CxHyOz) and char but also the light gas species distribution in volatiles, namely (CH4, CO, CO2, H2O and other light gases represented by C2H4). In the present work, the model prediction of volatile species was coupled with chemical kinetics for the gas species combustion as mentioned before. For the char conversion the oxidation reaction producing CO due to the small particle sizes and large temperature the oxidation reaction was considered first order with the pre-exponential constant A ¼ 0.005 kg/sm2Pa and activation energy 74,000 kJ/kmol. The thermal radiation is calculated using the discrete ordinates method. The radiative properties of the participating medium were modelled by the weighted-sum-of-grey gases model (SLW) [18] and the radiation properties of particles are calculated from approximations to the Mie theory [19]. The model considers distinct approximate functions to calculate the absorption and scattering coefficients of coal, char and ash particles.
3. Test cases The conditions considered for the simulations were those from the tests made at RWEn power's 0.5 MWth combustion test facility (CTF) [14]. As shown in Fig. 1 the CTF has a refractory lined combustion chamber composed by a block with 0.8 0.8 m inner cross-
2. Numerical model The model is based on the numerical solution of the Favreaveraged balance equations for mass of individual species, momentum, turbulence quantities and energy. The SIMPLER algorithm was used for velocity-pressure coupling and the standard k-ε model was used to model the turbulence. Combustion in the gas phase is calculated based on the Eddy Dissipation Model [15]/Finite Rate EDM/FR approach to describe
Table 1 Parameters for global reaction rates (*modelled as a long chain hydrocarbon with molecular weight of 120 kg/kmol). Fuel
A
CH4 C2H4 Tar* CO
5.01 2.42 1.98 2.24
1011 1010 1010 1012
Ea/R
a
b
c
Ref.
24,334 15,083 15,083 20,110
0.7 0.10 0.25 1.00
0.8 1.65 1.50 0.25
0 0 0 0.5
[16] [17] [17] [16]
A. Rebola, J.L.T. Azevedo / Applied Thermal Engineering 83 (2015) 1e7
3
Fig. 1. Schematic representation of the CTF [14].
section and 4 m long and a final convergent block with 2.18 m length and an end cross-section of 0.3 0.3 m. The installation is equipped with a 0.5 MWth (IFRF) aerodynamically air staged burner, a schematic representations of the burner is presented in Fig. 2. The CTF was originally designed for air-firing and was retrofitted to once-through oxyfuel system and the conditions considered in the present paper are those for a Russian coal fired with 3% O2 in the flue gases simulating dry recycling. The recirculation ratio, RR, is defined as:
RR ¼
mRFG mRFG þ mPFG
(2)
mRFG is the recirculated gas mass flow rate and mPFG is the flue gas mass flow rate. Table 2 shows the test conditions for the case of air firing and oxycoal combustion for the four recirculation ratios that were considered for the simulations. The temperatures of the primary and secondary stream were 70 and 270 C, respectively. The O2 concentration in the primary stream was maintained at 21% (V/V) for all recycle ratios and the primary stream velocity was maintained constant. Table 3 presents the composition and calorific
value of the Russian coal. The particle size distribution was approximated by a Rossin Rammeler distribution based on the diameters distributions of the Russian coal presented in Ref. [20] and shown in Table 4. The calculations were carried out using a staggered, two dimensional axisymmetric, non-uniform grid. The grid had a higher density near the burner region and in the vicinity of the centreline containing 126 120 cells and the computational domain extends over 7822 451 mm (axial radial direction). The radial dimension was obtained maintaining the cross section area of the axisymmetric mesh equal to the real furnace, and the axial dimension of the grid guarantees that the exit boundary is more than ten diameter away of the last cross area change. The boundary at the burner axis was treated as axisymmetric, and the downstream boundary was treated as outflow. At the burner exit the radial velocity component was taken as zero, the secondary stream tangential velocity was estimate based on the swirl number and the axial velocity was calculated from mass flow rate of the transport and secondary fluid presented in Table 2. The inlet turbulent kinetic energy, kin, was estimated assuming that the inlet turbulence intensity is 10%. The inlet values of the dissipation
Fig. 2. Schematic representation of the burner [14].
4
A. Rebola, J.L.T. Azevedo / Applied Thermal Engineering 83 (2015) 1e7
Table 2 Test conditions (*N2 mass flow). Recirculation ratio [%]
Coal [kg/h]
Primary fluid O2 [kg/h]
Air RR 65 RR 68 RR 72 RR 75
68.8 68.8 68.8 68.8 68.8
25 25 25 25 25
Secondary fluid CO2 [kg/h] 83 130 130 130 130
*
[%] mass
Moisture Volatile matter Ash Fixed carbon Carbon Hydrogen Nitrogen Sulphur Oxygen Gross calorific value [kJ/kg]
6.23 33.55 11.98 48.27 65.91 4.59 2.09 0.34 8.86 27,098
h h i i 2pklðTw T∞ Þ 4 G εsTw ð1 εÞG þ hðTin Tw Þ Aint ¼ lnðRext =Rint Þ (3) G represents the incident heat flux at the wall, Tw the interior wall temperature, T∞ the exterior wall temperature considered 298 K, h the convection heat transfer coefficient, Ain the interior wall area, l the control volume length near the wall, s the StefaneBoltzmann constant, k the wall thermal conductivity, ε the wall emissivity, Tin the temperature of the inside medium at the wall,
Diameters [mm]
[%] mass
300 212 150 106 75
99.7 99.2 97.6 91.6 77.8
141 129 130 132 134
CO2 [kg/h] 463 226 278 364 448
*
O2 [%] v/v 21 44 39 33 29
4. Numerical results and discussion
rate of turbulent kinetic energy were determined from: εin ¼ C kin/l. The mixing length, l, was taken as 5% of the inlet area square root the constant C was set equal to 0.09. Unfortunately the CTF wall temperature and emissivity are not available. To perform the simulations the internal wall temperature distribution of each case was estimated based on the energy balance at the wall, assuming that the heat flux thought the refractory wall is equal to the radiation flux absorbed by the wall at each section of the furnace. The furnace outside wall temperature was considered 298 K, the values 0.3 m and 2.74 W/(mK) was considered to the wall thickness and the thermal conductivity of the refractory respectively and for the wall emissivity the value of 0.85 was considered. The wall temperature distributions obtained by this approach for each case are presented below. We also perform simulations for the Air firing conditions with different wall properties to evaluate the influence of the wall boundary conditions on the results. On these simulations the wall temperature was calculated inside the code based on the local energy balance at the wall, expressed as:
Table 4 Coal particle size distribution (experimental values) [20].
21 21 21 21 21
O2 [kg/h]
and Rint and Rext the interior and exterior radius of the wall respectively. Simulations A and B were performed with a wall thermal conductivity of 3.54 W/mK and a wall emissivity of 0.85 and 0.5 respectively. Simulation C was performed with a wall emissivity of 0.85 and a lager thermal conductivity 4.2 W/mK.
Table 3 Proximate and ultimate analyses of coal as received and averaged [14]. Russian coal
O2 [%] v/v
Fig. 3 presents the stream lines and coal particle consumption rate, predicted by the model, for the five study cases s. The predicted temperature and oxygen volumetric fractions distributions are presented in Fig. 4. For all cases the flow, in the near burner region, is characterized by a short internal recirculation zone, attached to the quarl, and a large external recirculation zone. The model predicts ignitions, for the five cases, inside the internal recirculations zone, near the quarl. As it is showed by the low level of oxygen and the relatively high temperatures predicted in this zone. The higher temperature computed inside the internal recirculation zone for the Air firing conditions comparatively to the cases with recirculation could be explained by the low mass flow rate, for the same amount of oxygen, of the burner primary stream and the lower heat capacity of the N2 comparatively to CO2. As shown in Table 2 for the recirculation cases to keep the same oxygen flow rate and volumetric fraction in the primary stream, as the Air case, it is necessary to increase the burner primary mass flow rate. Further due to the reduction of the secondary flow rate for the lowers recirculation ratio, the primary to secondary momentum ratio increases and therefore the effect of the swirled secondary stream is smaller leading to a less intense internal and external recirculation zones. These changes, in the near burner zone flow, have a significant influence on the flame structure. For the Air firing conditions and the higher recirculation rates the particle consumption occurs essentially inside the external recirculation zone and in the shear layer between the main jet and the external recirculation zone. For the lowers recirculation ratio, the flow capacity to drag the particles into the external recirculation zone is reduced, as a consequence, the conversion of the coal particles occurs essentially in the shear layer between the main jet and the external recirculation zone. The structure of the flame approximates to a jet type flame for the lowers recirculation ratio. This is corroborated by the computed distributions of oxygen and temperature. For the lower recirculation ratio the oxygen level are very high in the external recirculation zone, as a consequence of the high levels of oxygen in the secondary stream and the lower penetration of the particles in the external recirculation zone. These cases present high temperatures and low levels of oxygen in the shear layer between the main jet and the external recirculation zone, where essentially occurs the combustion process. The RR 75 is the case with the oxygen level and mass flow rate of the secondary stream closer to the Air firing conditions. The flow, oxygen distributions and coal particles mass release computed for these two cases have more similarities comparing to the others cases. These results clearly indicate that, for combustion with
A. Rebola, J.L.T. Azevedo / Applied Thermal Engineering 83 (2015) 1e7
5
Fig. 3. Stream lines and particle consumption distribution. (Wall temperatures presented in Fig. 5, wall emissivity 0.85).
recirculated flue gases to obtain flames with the same structure as the Air firing conditions it is necessary to maintain the burner aerodynamic and have the appropriate level of oxygen in the burner stream. The computed temperatures increase with the reductions of the recirculation ratio due to the reduction of the mass flow rate of gases, for the same thermal input from coal, as the recirculation ration decreases. Despite the higher recirculation ratios present an oxygen distribution and flow structure more similar to the Air firing conditions, the RR 72 is the case with the computed temperature distribution closer to the Air case. The lower temperatures predicted for the RR 75, comparatively to the Air case, are explained by the higher heat capacity of CO2 comparatively to N2, these two cases have similar burner mass fluxes. For the RR 72 the difference of the heat capacity between CO2 and N2 is compensated with the reduction of the CO2 mass flow rate comparatively to the N2 of the Air firing conditions.
Fig. 5 presents the wall temperature distribution used in the simulations obtained based on the experimental incident wall flux and the comparison between the experimental results and the computed heat wall incident fluxes, for the five cases studies. It can be observed that in general there is good agreement between the computational results and the experimental data. Nevertheless the computed wall incident heat fluxes over predict the measured values for all cases. The RR 72 is the case with the computed wall incident closer to the Air firing conditions. These two cases have also a similar computed temperature distributions as shown in Fig. 4. Fig. 6 presents a comparison between the calculated incident heat fluxes obtained for Air, RR 65 and RR 75, considering the prescribed wall temperature of the Air firing condition, presented in Fig. 5. The difference between the incident heat fluxes, for the different cases, is much lower than the previous results shown in Fig. 5. These results show the influence of the wall temperature on
Fig. 4. Temperature and O2 wet volume distribution. (Wall temperatures presented in Fig. 5, wall emissivity 0.85).
6
A. Rebola, J.L.T. Azevedo / Applied Thermal Engineering 83 (2015) 1e7
mass flow rate of the RR 75 comparatively to the Air case. An explanation for the results presented in Fig. 6 is the higher contribution of the medium to radiation flux, due the higher levels of CO2, on the RR 75 comparatively to the Air case. Fig. 7 presents the computed wall temperature, incident and absorbed radiation heat fluxes of simulations A, B and C, as shows the comparison between the results of simulation A and B, a significant change of the wall emissivity had a marginal effect on the results. The incident heat flux is part from the emitted and reflected walls and from the participation medium gas and particles. A decrease of the wall emissivity causes a reduction of the heat flux emitted by the walls and an increase of the reflected flux, the results suggest that the wall effects cancels each others, as a consequence changes of the wall emissivity value have a small effect on the computed results. The reduction of the wall emissivity causes
Fig. 5. Prescribed wall temperature and predicted wall incident heat flux. (Wall emissivity 0.85).
the computed wall incident heat fluxes. The higher fluxes are obtained for the RR 65 firing conditions, as explained before, the lower flow rate of gases for the same heat input from the coal will provide higher temperature and as a consequence higher fluxes. For the RR 75 conditions the predicted wall fluxes are slightly higher than the case of Air firing on the early section of the CTF. The predicted temperatures, in the near burner zone, of the RR 75 are slight lowers than the Air case, a consequence of the higher calorific value of the CO2 comparatively to the N2 and the superior burner
Fig. 6. Wall incident radiation flux for Air, RR 65 and RR 75 with the prescribed temperature of the Air case presented in Fig. 5. (Wall emissivity 0.85).
Fig. 7. Predicted wall temperature, incident and absorbed radiation heat flux for different wall properties. (Simulations A, B and C).
A. Rebola, J.L.T. Azevedo / Applied Thermal Engineering 83 (2015) 1e7
just a small reduction on the computed wall temperature and the radiated heat flux absorbed by the walls in the early section of the CTF. On the other hand, changes of the wall thermal conductivity have a larger influence on the results, particularly on the computed wall incident heat flux. The comparison between the results of simulation A and C shows, as expected, an increase of the computed wall temperature with the reduction of the considered wall thermal conductivity. As already explained, the higher temperatures of the wall and consequently of the interior medium increase the wall incident flux. The results also show that, a small increase of the computed wall temperature, of approximately 3%, increases the computed incident heat flux in more than 10% and reduces the absorbed wall flux in about 7%. These results clearly show the importance of the wall temperature especially on the accuracy of the computed incident heat flux at the wall. The direct comparison of incident heat fluxes without analysing the effect of wall temperature may provide misleading conclusions. This is not the case for water cooled membrane walls in boilers where the thermal resistance of the wall is much smaller and the wall temperature much lower. However for hot refractory wall test furnaces where there is a high dependency between the internal wall temperature and the incident heat flux, care should be given to the interpretation of the results. 5. Conclusions This work presents a simulation of coal combustion using air and mixtures of CO2 and oxygen as comburent, giving special attentions to the radiations wall incident heat fluxes and the flames structure. The computational results are compared with experimental data obtained in a 500 kWth combustion test facility at RWEn. The numerical model considers an axisymmetric approximation for the flow and coal simulation based on a Lagrangian method. The calculations indicate that the flame structure is modified by the operating conditions, with the reduction of the recirculation ratio due to the reduction of the secondary flow rate, the effect of the swirled secondary stream becomes smaller and the structure of the flame approximates to a jet type flame. The computed results also show that oxy-flames with a level of oxygen of 29% volume on the secondary burner stream and with a similar momentum as the Air firing conditions present a flame structure similar to the Air case flame. With the appropriate wall temperature, the radiation heat fluxes predicted by the model are in agreement and have the same trends of the experimental data. The numerical result shows that flames with near computed temperature distribution presents similar computed wall incident heat radiation fluxes. The numerical results also show that the computed wall incident heat fluxes are much more influenced by the wall temperatures than by the recirculation ratio or the CO2 levels. Therefore the analysis of the radiation fluxes obtained in hot refractory wall test furnaces should be done with care to avoid misleading conclusions.
7
Acknowledgements ^ncia e a TecThis work was supported by Fundaç~ ao para a Cie nologia (FCT), through IDMEC, under LAETA Pest-OE/EME/LA0022 and the project BOFCom financially supported by the Research Fund for Coal and Steel under contract RFCReCTe2006e00010.
References [1] T. Wall, Y. Liu, C. Spero, L. Elliott, S. Khare, R. Rathnam, F. Zeenathal, B. Moghtaderi, B. Buhre, C. Sheng, R. Gupta, T. Yamada, K. Makino, J. Yu, An overview on oxyfuel coal combustion-state of the art research and technology development, Chem. Eng. Res. Des. 87 (2009) 1003e1016. [2] G. Scheffknecht, L. Al-Makhadmed, U. Schnell, J. Maier, Oxy-fuel coal combustion-A review of the current state-of-the-art, Int. J. Greenh. Gas Control 5 (2011) S16eS35. [3] L. Cheng, S.Z. Yong, A.F. Ghoniem, Oxy-fuel combustion of pulverized coal: characterization, fundamentals, stabilization and CFD modelling, Prog. Energy Combust. Sci. 38 (2012) 156e214. [4] A. Molina, C.R. Shaddix, Ignition and devolatilisation of pulverized bituminous coal particles during oxygen/carbon dioxide coal combustion, Proc. Combust. Inst. 31 (2007) 1905e1912. [5] J. Brix, P.A. Jensen, A.D. Jensen, Coal devolatilisation and char conversion under suspension fired conditions in O2/N2 and O2/CO2 atmospheres, Fuel 89 (2010) 3373e3380. [6] R.K. Rathnam, L.K. Elliot, T.F. Wall, Y. Liu, B. Moghtaderi, Differences in reactivity of pulverised coal in air (O2/N2) and oxy-fuel (O2/CO2) conditions, Fuel Process. Technol. 90 (2009) 797e802. [7] K. Anderson, R. Johansson, S. Hj€ artstam, F. Johnsson, B. Leckner, Radiation intensity of lignite-fired oxy-fuel flames, Exp. Therm. Fluid Sci. 33 (2008) 67e76. [8] Y. Tan, E. Croiset, M.A. Douglas, K.V. Thambimuthu, Combustion characteristics of coal in a mixture of oxygen and recycled flue gas, Fuel 85 (2006) 507e512. [9] S.P. Khare, T.F. Wall, A.Z. Faruda, Y. Liu, B. Moghtaderi, R.P. Gupta, Factors influencing the ignition of flames from air-fired swirl pf burners retrofitted to oxy-fuel, Fuel 67 (2008) 1042e1049. [10] S. Leiser, Numerical Simulation of Oxy-fuel Combustion (Ph.D. thesis), University of Stuttgart, Germany, Shaker-Verlag Aachen, 2011, ISBN 978-3-83229722-0. [11] D. Toporov, P. Bocian, P. Heil, A. Kellermann, H. Stadler, S. Tschunko, € rster, R. Kneer, Detailed investigation of a pulverized fuel swirl flame in M. Fo CO2/O2 atmospheres, Combust. Flame 155 (2008) 605e618. € rster, R. Kneer, Development [12] P. Heil, D. Toporov, H. Stadler, S. Tschunko, M. Fo of an oxycoal swirl burner operating at low O2 concentrations, Fuel 88 (2009) 1269e1274. €ster, R. Kneer, Experimental and [13] M. Habermehl, J. Erfurth, D. Toporov, M. Fo numerical investigation on a swirl oxycoal flame, Appl. Therm. Eng. 49 (2012) 161e169. [14] J.P. Smart, P. O'Nions, G.S. Riley, Radiation and convective heat transfer and, burnout in oxy-coal combustion, Fuel 89 (2010) 2468e2476. [15] B.F. Magnussen, B.H. Hjertager, On mathematical modeling of turbulent combustion with special emphasis on soot formation and combustion, Proc. Combust. Inst. 16 (1977) 719e729. [16] F.L. Dryer, I. Glassman, High-temperature oxidation of CO and CH4, Proc. Combust. Inst. 14 (1973) 987e1003. [17] C.K. Westbrook, F.L. Dryer, Simplified reaction mechanisms for the oxidation of hydrocarbon fuels in flames, Combust. Sci. Technol. 27 (1981) 31e43. [18] M.K. Denison, A Spectral Line-based Weighted-sum-of-gray-gases Model for Arbitrary RTE Solvers (PhD thesis), Bringham Young University, United States of America, 1994. [19] W.A. Fiveland, W.J. Oberjohn, D.K. Cornelius, A Numerical Model for Predicting Furnace Performance in Axisymmetric Geometries, Department of Energy, United States of America, 1985. Report DOE/PC/40265e9. [20] J.P. Smart, R. Patel, G.S. Riley, Oxy-fuel combustion of coal and biomass, the effect on radiative and convective heat transfer and burnout, Combust. Flame 175 (2010) 2230e2240.